Compact, high-effectiveness, gas-to-gas compound recuperator with liquid intermediary

ABSTRACT

A liquid-loop compound recuperator is disclosed for high-ε heat exchange between a first shell-side fluid stream and a second shell-side fluid stream of similar thermal capacity rates (W/K). The compound recuperator is comprised of at least two fluid-to-liquid (FL) recuperator modules for transfer of heat from a shell-side fluid, usually a gas, to an intermediary tube-side heat transfer liquid (HTL). Each FL module includes a plurality of thermally isolated, serially connected, adjacent exchanger cores inside a pressure vessel. The cores are rows of finned tubes for cross-flow transfer of heat, and they are arranged in series to effectively achieve counterflow exchange between the HTL and the shell-side stream. The HTL may be water, an organic liquid, a molten alloy, or a molten salt.

CROSS-REFERENCE TO RELATED APPLICATIONS

This application claims the benefit of U.S. application No. 61/016,247filed Dec. 21, 2007, and the benefit of U.S. application No. 61/034,148filed Mar. 5, 2008, each of which is incorporated herein by referencefor all purposes.

FIELD OF THE INVENTION

The field of this invention is heat exchangers, and more particularly,compact, gas-to-gas recuperation at high effectiveness for clean gasesof similar heat capacity rates using compound recuperators with liquidintermediary.

BACKGROUND OF THE INVENTION

Gas-to-gas recuperation with both high thermal effectiveness andorder-of-magnitude improvement in cost effectiveness is critical toaddressing global energy needs, as shown in at least two co-pendingpatent applications. From a manufacturing perspective, the challengesarise from the fact that it is not practical to produce heat exchangerswith closely-spaced fins on both the inside and outside of tubes, andalternative approaches thus far have had limited success.

An enormous number of heat exchangers have been well optimized fornumerous purposes over the past four decades. However, most have notbeen directed at high thermal effectiveness ε for cases where the heatcapacity rates in the two streams are similar. The heat capacity rate Wof a stream is given by GC_(P) (its SI units are W/K), where G is themass flow rate (kg/s) and C_(P) is the specific heat (J/kg-K). By thestandard definition of ε (the ratio of heat transferred to thetheoretical limit), high ε is most easily achieved when W_(min) (that ofthe weaker stream) is much less than W_(max) (that of the strongerstream). However, exergy destruction can be minimized only if W_(min) isclose to W_(max). The terms “recuperator” and “regenerator” have usuallyimplied the streams have similar W's, and that will be the usage andregime of primary focus in this invention. However, the streams need notbe in the same state—one may be liquid while one is a gas.

Common examples of cost-effective heat exchangers with high exergy lossinclude automobile radiators and air-conditioning condensers. In theautomobile radiator, for example, the warmed air leaves at a temperaturemuch below that at which the hot water enters. Thus, most of the water'sexergy (energy availability) has been destroyed, irrespective ofprecisely how one chooses to define it. Other examples of cost-effectivecompact exchangers for unrelated purposes include micro-channel,compact, fluid cooling systems, as seen for example in U.S. Pat. No.6,907,921.

The subset of fluid heat exchangers directed at high ε have mostlyaddressed one of the following cases: condensing-vapor-to-liquid,condensing-vapor-to-gas, boiling-liquid-to-liquid,boiling-liquid-to-gas, liquid-to-gas, or liquid-to-liquid. In all ofthese cases, the fluid thermal conductivities, k_(t), W/m-K, are fairlylarge on at least one side (generally over 0.2 W/m-K), or phase changeis present to drive small-scale turbulence on one side. A commongas-to-gas exchange application is in steam power-plant superheaters.However, the steam here has high thermal conductivity and rather highdensity (for example, 0.067 W/m-K and 40 kg/m³ at 10 MPa, 650 K).Moreover, high ε there is not an objective, as the flue gas will be usedsubsequently for boiling. Gas-to-gas exchange is also sometimes seen inair preheaters in steam power plants. Here, moderately high ε may beseen, though usually the minimum flue-gas exhaust temperature is ˜400 Kto limit corrosion from acid condensation, and this limits ε of theserecuperators:

Achieving high ε in gas-to-gas exchange with low pumping power has beenchallenging because volumetric specific heats are much lower than seenin liquids and thermal conductivities are usually low. Challenges arealso seen in achieving high ε in recuperators for organic liquids justabove their pour point where viscosity is quite high.

Doty, in U.S. Pat. No. 4,676,305, disclosed a compact method ofachieving highly effective recuperation with low pressure drop for gasesof similar W's. However, this microtube recuperator has not yet beenshown to be commercially competitive with the brazed plate-fin type, inwide usage in recuperated open Brayton cycles in the 30-250 kW range andoccasionally up to 25 MW. See, for example, the microturbines availablefrom Capstone Turbines Corporation, of Chatsworth, Calif. These too havelimited cost effectiveness and limitations in accommodating applicationswhere there are large pressure differences (greater than ˜0.7 MPa)between the two streams at high temperatures (above ˜750 K).

Optimized, compact high-ε gas-to-gas recuperators require low flowvelocities (several percent of the sonic velocity), total flow-pathexchange lengths in the range of 0.1 to 2 m, and passage hydraulicdiameters of 0.5 to 8 mm, with the larger diameters corresponding topressures near 0.1 MPa and the smaller sizes corresponding to pressuresabove 0.5 MPa. They have also required the use of construction materialshaving fairly low thermal conductivity, though that is not required inthis invention.

An alternative to paralleling tens of thousands of microtubes that hasseen rather little usage but appears to be the most competitive for somecompact recuperation applications is the rotating honeycomb regenerator,as used in some turbine engines where system mass is critical. Oda et alin U.S. Pat. No. 4,304,585 disclose an early ceramic design.Regenerators have seen very little usage largely because of thedifficulties in obtaining adequate isolation between the high-pressureand low-pressure streams and because of the shedding of ceramicparticles, leading to turbine abrasion.

Ceramic is usually selected for honeycomb regenerators in recuperatedaero-turbine applications because of the need for oxidation resistanceat high temperatures and the advantage of low thermal conductivity inthe flow direction. Rotating ceramic honeycomb regenerators havedemonstrated ε above 98% while the brazed plate-fin recuperators seldomachieve more than 87% ε, primarily because of cost and mass optimizationconstraints. The honeycomb regenerators can be an order of magnitudemore compact and an order of magnitude less costly for a given exchangepower and ε than plate-fin microturbine recuperators—which are an orderof magnitude more compact than the gas-to-gas exchangers seen in mostcurrent chemical engineering and power generation applications.

Oxidation resistance is irrelevant in some applications, and therehoneycomb regenerators can be made at lower cost and with much higherreliability from a low-conductivity alloy honeycomb, such as siliconbronze, stainless steel, or some magnesium or aluminum alloys. Thethermal conductivity of silicon-nickel-bronze can be below 40 W/m-K, and120 W/m-K is sufficiently low except perhaps for the most compactapplications. For example, a magnesium alloy with thermal conductivity˜90 W/m-K has been used experimentally in a helicopter turboshaftengine. Titanium alloys would be better, and their relative cost shoulddecrease over the next decade. The much higher thermal stress toleranceof metals compared to ceramics is extremely beneficial with respect todurability, as thermal stress is a primary factor limiting ceramicregenerator design and contributing to shedding of particles fromceramic regenerators.

Regenerator cost for a given performance is typically near minimum whenpore diameters are about 0.7 mm for many mobile gas-gas exchangeapplications. The relevant design theory, well understood for more thanthree decades, has recently been reviewed and updated by David G Wilsonin “Design and Performance of a High-Temperature Regenerator Having VeryHigh Effectiveness, Low Leakage and Negligible Seal Wear”, paper GT2006-90096, Turbo-Expo 2006. The use of a metal for the honeycomb,possibly with the innovations in Wilson's U.S. Pat. No. 5,259,444, maypermit a satisfactory solution of the sealing and wear problems inlarger recuperators where the pressure difference between the twostreams is small.

However, the rotating honeycomb regenerator still has substantiallimitations, either where there are substantial pressure differencesbetween the two streams, or where the size is small (below ˜100 kW), orwhere the lower-pressure stream is above ˜0.4 MPa. This last conditionleads to greater difficulties in limiting leakage and carry over, and itleads to unreasonably low porosity requirements (or high solidity) inthe honeycomb (for sufficient thermal storage). High solidityexacerbates axial thermal conduction losses and makes the regeneratormore massive and perhaps more prone to stress-related failure. When twoor more of the above conditions are present simultaneously, thehoneycomb suffers markedly.

High-ε recuperators are essential in many cryogenic processes. A commonand extremely effective design in cryocoolers uses micro-multi-port(MMP) tubing with one of the gases flowing in one direction through someof the “ports” (passages) and the other flowing in the oppositedirection through the other ports. The viscous losses in very longlengths (4-20 m) of microtubes (under 1 mm ID, inside diameter) areoften fully acceptable for gases at very high pressures (over 1 MPa) andlow temperatures (below 140 K). Many cryogenic recuperators operate atsuch conditions, where outstanding counterflow recuperators can be madefrom MMP tubing or a similar construction. For many cryogenicapplications outside the above conditions, the novel compoundrecuperator presented herein will be superior.

The basis for the innovation presented herein begins by learning fromthe highly developed liquid-to-gas exchangers best exemplified inair-conditioning (AC) condensers and automobile radiators. To achievethe high ε sometimes needed in gas-to-liquid recuperation, it is simplynecessary to arrange 5 to 30 of such exchangers in series, with theliquid flowing serially from the first to the last and the gas alsoflowing serially, but from the last to the first. Such a counterflowexchanger can be an order of magnitude less massive and less costly thanconventional shell-and-tube gas-to-liquid counterflow exchangers ofsimilar flow rates, pressure drops, and ε.

The dry-air condensers ubiquitous in AC condensers have been extremelywell optimized by numerous air-conditioning companies over the past fourdecades. For example, “80-ton” (280 kW of cooling) air conditioners arewidely produced. The air-flow passage lengths in these condensers areoften under 3 cm per row of tubes; and air-passages, though perhapsseveral centimeters wide, typically have thicknesses of ˜1.5 mm. Thiscorresponds to a hydraulic diameter of ˜3 mm for the air flows, whichinterestingly is that predicted to be optimum at 0.1 MPa by thealternative analysis presented by Doty in U.S. Pat. No. 4,676,305. Thecondenser in such a unit typically rejects about 350 kW_(T) at a δT (dryair) of about 10° C. Some large commercial freezer systems utilizerefrigerant R744, CO₂, where condenser pressures can exceed 6 MPa, soclearly high tube-side pressures can be accommodated by cross-finnedtubes produced by automated manufacturing processes as used in ACcondenser cores. These exchanger cores are usually intended for use withtwo-phase flow tube-side over a significant portion of their length.However, predominately single-phase tube-side liquid flow, as seen forexample in U.S. Pat. No. 3,922,880 in a design for use in an AC unit,can also be very cost effective.

In U.S. Pat. No. 4,831,844 Kadle discloses that for condensing two-phasetube-side flow, substantial improvement is obtained by a step-downapproach in which the tube-side vapor flow begins in two parallel tubesand then combines to a single tube about two-thirds of the way throughthe condensing process. Several advantages are noted for many ACapplications, but the drawings therein also appear to show interleavedtube-side flow between parallel rows of finned tubes. Both step-down andinterleaved tube-side flow would generally be disadvantageous forsingle-phase tube-side liquid flow in high-ε exchange, as addressedherein; but with such patterns avoided, common AC condenser cores may beutilized for high-ε recuperation.

Another common approach to improving tube-side heat transfer with alow-velocity liquid of low k_(t) is to use MMP tubing for theliquid-phase flow, as discussed by Guzowski et al (IMechE 1999,C543/083) and Guntly et al (U.S. Pat. No. 4,998,580). A simpler methodis to insert turbulators, such as open-pitch coil springs inside thetubes. This can be quite beneficial with single-phase tube-side flow ofliquids under certain conditions.

The solution for order-of-magnitude improvement (compared toshell-and-tube exchangers) in cost-effective high-s recuperation betweenclean gases shell-side at moderate temperatures and liquids tube-side isto simply use a series arrangement of several AC condenser cores (ofproper design), serially connected inside a pressure vessel. Asseemingly obvious and advantageous as the above approach is for high-εgas-liquid exchange, it does not appear to have been practiced assuch—liquid-only tube-side flow through a series of thermally isolatedcores. Related exchangers, in which the shell-side gas goes cross-wiseback and forth several times over the length of cross-flow tubes, arecommonplace; and often the tubes have fins (though usually spaced 3 to15 mm). However, the above differences are of enormous importance withrespect to manufacturing, compactness, and cost effectiveness.

Single- and multi-row cores similar to what are suitable for a componentin the instant invention are produced by Armstrong under the productname Duralite™ Plate Fin Coils. But apparently the value of thermallyisolated serially connected cores inside a pressure vessel has notpreviously been appreciated as optimum to achieve high-ε.

Perhaps series arrangements of thermally isolated cores similar to thoseused in AC condensers have not been considered for high-ε gas-liquidexchange because most large applications also require dealing withmoisture, acids, and particulates in the gas stream. For many suchcases, available shell-and-tube exchangers, developed primarily forcondensing shell-side steam, with typical tube diameters of 12 to 50 mmand shell-side fins usually spaced ˜6 mm, may be the best option,especially when the gas pressure is below 0.12 MPa and high ε is notdesired.

The heat pipe is in some sense related to the compound exchangerdisclosed herein, as it too uses an intermediary fluid. However, theheat pipe uses a self-pumped two-phase fluid tube-side, and it is poorlysuited to gas-gas recuperation. So the relationship to the heat pipe istenuous at best. A complex, finned, device cooler shown in U.S. Pat. No.7,296,619 may incorporate heat pipes, though that document tries todistort and confuse the standard meaning of “heat pipe”. Regeneratorsare also somewhat related, as they utilize an intermediary, but there itis a solid.

The standard air conditioner is most closely related to the inventivecompound recuperator, as it too provides heat transfer between two gasesusing a fluid intermediary. There, however, the large majority of theheat transfer in each exchanger includes phase change, and a veryenergy-intensive vapor pump is required. It is possible that someair-to-air recuperators for heat recovery in buildings have utilizedproprietary concepts somewhat related to those presented herein, butapparently all such have relied upon phase change in the fluidintermediary for most of the heat transfer, and there is no evidencethat they have achieved high ε.

Tube-side phase change has previously been desired because it greatlyincreases tube-side heat transfer coefficient, h_(t), W/m²-K, and thusgenerally allows significant reduction in exchanger size. However, phasechange is not desired in the instant invention, as it makes minimizationof irreversibilities impractical (because it requires a very largenumber of intermediary loops). The instant invention allows for enormousreduction in exchanger size without phase change, and thus it alsoreadily permits high ε. Not surprisingly, commonly used “refrigerants”are the worst type of fluids that can be imagined for the applicationsenvisaged by the instant invention.

It is noteworthy that the chemical engineering process simulationsoftware we have evaluated is not capable of handling the case where atube-side liquid stream is being heated by gas in a cross-flowfinned-tube exchanger, as seen in the instant invention.

Two co-pending patent applications disclose enormous, emergingapplications for high-ε low-cost recuperation between clean gases wheregood solutions are not currently available: (A) where the hot gas streamenters above 550 K and at more than 0.2 MPa, especially if the pressuredifference between the streams exceeds 1 MPa, (B) where some liquidcondensation or frosting can be expected in one or both of the gasstreams, and (C) where both gases are at pressures below 1 MPa, thepressure difference exceeds 0.1 MPa, the temperatures are above 90 K,and cross-contamination must be avoided. There also appears to be anenormous, emerging application for high-ε low-cost recuperation betweenviscous organic liquids. The invention presented herein addresses theseand many other situations most optimally.

The instant invention is, in practice, usually implemented as a minimumof two separate modules with one or more liquid intermediary loopsbetween them. Naturally, each independent module is usable as afluid-to-liquid recuperator, where the shell-side fluid is usually a gasbut may be a viscous liquid of low thermal conductivity.

RELEVANT ART

-   1. M M Guzowski, F F Kraft, H R McCarbery, J C Noveskey, “Alloy and    Process Effects on Brazed Automotive Condenser Tubing”,    http://www.ent.ohiou.edu/˜kraft/VTMS4paper.pdf, presented at IMechE    1999, C543/083.-   2. F D Doty, G Hosford, J B Spitzmesser, and J D Jones, “The    Micro-Tube Strip Heat Exchanger”, Heat Transfer Engr., 12, 3, 31-41,    1991.-   3. D G Wilson and J Ballou, “Design and Performance of a    High-Temperature Regenerator Having Very High Effectiveness, Low    Leakage and Negligible Seal Wear”, paper GT 2006-90096, Turbo-Expo    2006, Barcelona.-   4. Trane Product Literature, “Installation, Operation, Maintenance:    Series R”,    http://www.trane.com/webcache/rf/rotary%20liquid%20chillers%20(rlc)/service/rtaa-svx01a-en_(—)09012005.pdf,    RTAA-SVX01A-EN, 2005.-   5. F P Incropera and D P Dewitt, “Introduction to Heat Transfer”,    Wiley, NY, 2002.-   6. R K Shah, A D Kraus, D Metzger, “Compact Heat Exchangers”,    Hemisphere Pub., NY, 1990.-   7. L R Rudnick, “Synthetics, Mineral Oils, and Bio-based Lubricants:    Chemistry and Technology”, CRC, Boca Raton, 2006.-   8. K Weissermel, H J Arpe, Industrial Organic Chemistry, 4th ed.,    Wiley, 2003.-   9. C H Bartholomew and R J Farrauto, Industrial Catalytic Processes,    Wiley, 2006.-   10. E Prabhu, “Solar Trough Organic Rankine Electricity System    (STORES)”, NREL/SR-550-39433,    http://www.nrel.gov/docs/fy06osti/39433.pdf, 2006.-   11. DESIGN II for Windows, Version 9.4, 2007, by WinSim Inc.,    documentation available from    http://www.lulu.com/includes/download.php? fCID=390777&fMID=810115.-   12. Armstrong Duralite™ Plate Fin Coils, product information,    Granby, Quebec, 2008,    http://www.armstronginternational.com/files/common/allproductscatalog/platefincoils.pdf

U.S. PATENT DOCUMENTS 3,922,880 December 1975 Morris  62/498 3,994,337November 1976 Asselman et al 165/119 4,304,585 December 1981 Oda et al 65/43 4,645,700 February 1987 Matsuhisa et al 428/116 4,676,305 June1987 Doty 165/158 4,831,844 May 1989 Kadle  62/507 5,259,444 September1993 Wilson 165/8 5,435,154 July 1995 Nishiguchi  62/476 6,907,921 June2005 Insley 165/170 6,957,689 October 2005 Ambros et al 165/41 7,225,621June 2006 Zimron et al  60/651 7,296,619 November 2007 Hegde 165/104.33

U.S. PATENT Application Publication US 2006/0211777 September 2006Severinsky

SUMMARY OF THE INVENTION

A liquid-loop compound recuperator is disclosed for high-s heat exchangebetween a first shell-side fluid stream and a second shell-side fluidstream of similar thermal capacity rates (W/K). The compound recuperatoris comprised of at least two fluid-to-liquid (FL) recuperator modulesfor transfer of heat from a shell-side fluid, usually a gas, to anintermediary tube-side heat transfer liquid (HTL). Each FL moduleincludes a plurality of thermally isolated, serially connected, adjacentexchanger cores inside a pressure vessel. The cores are rows of finnedtubes for cross-flow transfer of heat, and they are arranged in seriesto effectively achieve counterflow exchange between the HTL and theshell-side stream. The HTL may be water, an organic liquid, a moltenalloy, or a molten salt. Alumina-dispersion-strengthened-metal fins,superalloy tubes, and a lead-bismuth-tin alloy HTL may be used for hightemperatures. Cumene may be used as the HTL in cryogenic applications.

BRIEF DESCRIPTION OF THE DRAWINGS

FIG. 1 illustrates schematically a multi-stage, liquid-loop, compoundrecuperator.

FIG. 2 illustrates the preferred liquid routing for a portion of acompound exchanger.

FIG. 3 is a perspective, cut-away view of a typical fluid-liquidexchanger module.

FIG. 4 illustrates a typical, single-row, finned-tube core.

FIG. 5 illustrates a serpentine pattern in a finned-tube core.

FIG. 6 illustrates five thermally isolated series tubes.

FIG. 7 illustrates a radial-flow version of an FL module.

DETAILED DESCRIPTION OF THE PREFERRED EMBODIMENT

FIG. 1 illustrates a 4×3 array of 12 liquid-gas cross-flow exchangercores with 2 liquid pumps and two different heat transfer liquids as anexample of a method of achieving high-c recuperation between twoisolated fluids of low thermal conductivity, gas-1 and gas-2, identifiedin the figure using hollow lines. These fluids have mean thermalconductivity less than 0.4 W/m-K (that of H₂ at ˜720 K) and will usuallybe gases with k_(t) less than 0.06 W/m-K. Thus, for improved clarity,they are generally referred to as gases herein, though applicationswhere these fluid streams would be viscous organic liquids are seen in aco-pending patent application. Both gas-1 and gas-2 are shell-side,sometimes also called “fin-side”. In this example, gas-1 is the hotsource stream, and gas-2 is the cold stream being heated to nearly theentry temperature of gas-1. Often, the hotter gas will be at lowerpressure than the cooler gas, but the reverse relationship is alsopossible.

In the example of FIG. 1, there are four sets of exchangers (A, B, C,D). The heat transfer liquids (HTLs) are identified in the figure withheavy solid lines. Here, each is directed serially through threecross-flow exchanger cores for each gas stream. The HTLs are alltube-side.

In this example, gas-1 enters 1 fin-side into exchanger labeled D1 at760 K and exits 2 fin-side from exchanger B3 at 400 K. Gas-2 enters 3fin-side into exchanger labeled A1 at 320 K and exits 4 fin-side fromexchanger C3 at perhaps 680 K. For such temperatures with similar Ws, εwould be about 78% by the standard definition.

Here, each gas stream passes through 6 cross-flow exchanger cores, threeon each side of each liquid loop. In practice, this will often be aminimum number, though it also depends on how one defines a cross-flowexchanger core. For example, a typical AC condenser “core” contains 2,3, or 4 rows of finned tubes, often connected serially. Hence, atypical, 3-row, serial “AC core” could perform the three serialexchanges as shown in FIG. 1 for each side of each loop. Herein, 3 rowsof thermally isolated finned tubes, serially connected, is considered tobe three cross-flow exchanger cores in series. For the rows to beconsidered thermally isolated, it is required that the fin metal not becontinuous from one row to the next (at least on most of the fins) andthat the tube flow pattern not be interleaved—that is, that the tubesnot return back to a first row after leaving that row and going to asecond row. From a functional perspective, the rows may be considered tobe thermally isolated if the thermal conduction of the metal betweenadjacent rows is less than twice the thermal conduction of the fluids(the sum of the shell-side and tube-side) between the rows.

For variety in presentation, a thermally isolated, serially connected,cross-flow exchanger core may be referred to as a “finned tube row”. Thecomplete serial group of tube rows in a single HTL loop for one of thegases will be referred to as a “core set”. The core sets will be insidea pressure vessel to contain the shell-side pressure, and often all thesets associated with the first gas stream would be inside one pressurevessel, and those associated with the second gas stream would be insidea second pressure vessel. For example, sets B and D of FIG. 1 wouldnormally be inside one pressure vessel and sets A and C would normallybe inside a second pressure vessel. The pressure vessel with the coresit contains may be referred to as a fluid-to-liquid (FL) recuperatormodule or a gas-to-liquid (GL) recuperator module, as the shell-sidefluid will usually be a gas.

The combination of two FL recuperator modules coupled with anintermediary HTL may be referred to as a liquid-loop recuperator or acompound recuperator. At least one liquid pump 5 and surge tank 6 arealso required for each compound recuperator. FIG. 1 illustrates adual-loop compound recuperator.

For minimization of 6T-related irreversibilities, the thermal capacityrate W_(L)=G_(L)C_(PL) of the HTL through a core set in a compoundrecuperator should be close to the geometric mean Ws of the thermalcapacity rates W₁ and W₂ for the two shell-side gas streams, G₁C_(P1),and G₂C_(P2),

W _(L) ˜W _(S)=(W ₁ W ₂)^(0.5).  [1]

Moreover, the ratio W₁/W₂ should be fairly close to 1, though thecompound recuperator will also be advantageous for other conditions.Normally, W_(L) would be between 0.7W_(S) and 1.4W_(S). Of course, G_(L)is proportional to npvd², where n is the number of parallel tubes in acore, ρ is the fluid density, ν is the flow velocity, and d is the tubeinside diameter (mm).

The practical ε limit (for similar W's) is essentially determined by thetotal number of rows, n_(r), of isolated, serially connected, finnedtubes (or cores) and the “number of heat transfer units”, NTU, where

NTU=h _(tS) A _(X) /W _(S),  [2]

where A_(X) is the heat transfer surface area. The ε suggested for FIG.1 is probably above a cost-effective limit for just 12 total cores withliquid intermediaries, though it is certainly possible. On the otherhand, with 16 cores per set, four sets, and two liquid loops, apractical limit of about 94% would be expected. The same practical limitwould be expected with a single liquid loop and 32 cores per set. Such adesign would be preferred when the temperature difference between thehot source gas and the cold source gas is rather small, as this requiresonly one liquid pump. Having multiple loops, as shown in the multi-stagecompound recuperator of FIG. 1, allows for the use of different HTLs indifferent temperature ranges, which allows for improved performance inlarge recuperators operating over a large temperature range.

FIG. 1, though drawn acceptably by diagramming conventions and chosenfor its clarity, does not convey flow details that improve maximumpractical effectiveness per tube row. The fluid routing shown in FIG. 2better conveys liquid routing details that significantly improveeffectiveness per stage. There, the liquid enters each row from the sameside relative to the shell-side flow, which is always distributed acrossthe face of the tube rows, as indicated by using parallel gas-flowarrows. The object is to make the direction of the thermal gradientalong each row the same and maintain a fairly uniform change in the gastemperature per row across the face of each core. Such a tube-side flowpattern is uncommon in AC condenser cores, as ε there is not soimportant.

Suitable finned-tube AC condenser or evaporator cores, though generallywithout the most optimum tube-side flow routing, are readily availablefor efficient heat transfers at power levels from about a hundred wattsto tens of kilowatts, and heat transfers of hundreds of megawatts can behandled just as cost effectively by paralleling tens of thousands ofsuitable AC cores. AC condenser cores are available at low cost becauseefficient production methods have been so highly optimized from thecompetitive pressures of high-volume manufacturing over the past fourdecades. Accommodating high shell-side pressures is straightforward—onesimply places the assembly in a large pressure vessel with suitablebaffles, as seen in U.S. Pat. No. 4,676,305, for example, and addressedlater in more detail. Although AC condenser cores are usually intendedfor operation near 310 K, they are sometimes constructed using coppertubing with aluminum or copper fins brazed on using a filler materialhaving liquidus near 870 K. It is also not too uncommon to use 90Cu-10Nialloy C706 for the tubing with copper fins. In larger sizes these corestypically use tubing of 9 to 13 mm diameter, and the fin pitch(center-to-center spacing) is often under 2 mm. Fin length in thedirection of air flow is typically ˜25 mm per row, though sometimes upto 80 mm per row. Fin pitch in the FL recuperator up to 8 mm may bedesired if the shell-side fluid is a very viscous liquid, such as an oiljust above its pour point.

Only minor modifications of available cores are needed to permitoperation to about 700 K at limited pressures with non-oxidizing cleangases. Moreover, operation to 900 K in many non-oxidizing conditions ispossible by simply changing to an alumina-dispersion-strengthened coppersuch as C15720 (0.4% Al₂O₃, bal Cu) for the fins and to the common70Cu-30Ni alloy C715 for the tubing (˜70 MPa yield strength at ˜900 Kfor C715, compared to ˜750 K for alloy C706).

The benefits of this approach may not be immediately clear to thoseaccustomed to evaluating heat transfer primarily on the basis of surfacearea, as (A) changing from a conventional shell-and-tube exchanger to atypical AC condenser core may increase the shell-side surface area pervolume by only a factor of 5 to 10 (from ˜200 m²/m³ to 1000 or perhapseven 2000 m²/m³), (B) the tube-side “compactness ratio” may decrease bya factor of 2 or more, and (C) the heat has to be transferred twice.What may be overlooked is that the shell-side heat transfer coefficient,h_(t), W/m²-K, will also typically increase by a factor of 5 to 10because the passage thicknesses are decreased, so the shell-side totalbenefit can be a factor of 25 to 100. By selection of an optimum HTL andflow velocity, the tube-side h_(t) can easily be made over 30 times(possibly even more than 200 times) that of most gases—usually withoutadding tube-side turbulators. Hence, the novel compound-exchanger canpermit an order of magnitude improvement in compactness compared toshell-and-tube exchangers for gas-gas exchange (for comparable powers,flow rates, ε, and pumping losses) even though the heat has to betransferred twice.

The increased complexity associated with compound recuperators using aliquid intermediary would not be justified below some size threshold.That cutoff is dependent on many variables, including desired ε,temperatures, gas compositions, the importance of mass reduction,cleanliness of the gas streams, and gas pressure differences. It willalso depend on availability of appropriate finned-tube cores for therelevant conditions, a factor that is likely to change markedly overtime. Even today, it seems that compound recuperators would be preferredfor many cases with non-oxidizing gases if ε above 70% is desired at (A)temperatures below 700 K, (B) exchange power levels above 20 kW, and (C)mean gas pressures above 0.05 MPa. Suitable cores for competitivecompound exchangers for a much wider range of conditions should becomeavailable.

A few more comments are useful to help elucidate the value and henceinventiveness of the instant invention. The shell-side thermal specificconductance, W/kgK, under the typical shell-side flow conditions(largely laminar) will be inverse with the square of the pitch; but themass will be nearly independent of the pitch for a given core volume.Clearly, as materials become steadily more expensive, there will be astrong incentive to minimize the pitch to permit ever higher heattransfer per exchanger mass. Of course, the shell-side pressure dropwill increase inversely with the pitch for a given flow velocity.However, most applications will be with shell-side pressure well above0.2 MPa, and the shell-side pumping power losses will often be inversewith the square of the gas density. Hence, smaller fin pitch than isgenerally seen in AC condenser cores will often be optimum, as long asthe shell-side flow-section area A_(S) (the frontal section area, notA_(X)) is kept large and the flow path length is kept short, asdiscussed later in more detail.

Minimum channel thickness in most prior-art, compact, high-effectivenessexchangers is ultimately limited by the need to establish highly uniformflow. Hence, manufacturing tolerances limit minimum spacing. Channelthickness tolerance is not as critical in the instant invention becauseflow mixing can readily occur between successive thermally isolatedcores.

Current AC condenser practice (˜2 mm fin pitch) is probably close tooptimum in the compound exchanger for mean gas pressures of ˜0.3 MPa,mean k_(t) of ˜0.04 W/m-K, and ε of 75-90%. The fin pitch can often bereduced (for further reductions in exchanger mass and cost) at highergas pressures or low temperatures. However, there are limits, as the finthickness must be sufficient to provide the needed thermal conductionand stiffness, and corrosion lifetime may be an issue in some cases. Astrong advantage of the liquid-loop compound exchanger compared toregenerators is that the gas-passage thickness of the high-pressure gasstream may readily be made much less than that for the low-pressure gasstream, as desired for maximum performance.

It is not uncommon in AC condensers for the tubes to have internalfeatures such as ribs, fins, or undulations to increase h_(tL), thoughthis adds to tubing cost and increases stress concentrations. Suchsurface enhancement is mostly beneficial in the initial portion of acondenser where no condensation is occurring (the tube-side vapor isstill superheated) and in the final portion (subcooling) where theliquid velocities are very low. Surface enhancements are much lessbeneficial in the compound recuperator because the tube-side flow isliquid-only, of essentially constant velocity, which may be betteroptimized.

Heat-Transfer Liquids (HTLs). The primary requirements in the HTL arechemical stability at the relevant conditions, low viscosity, low vaporpressure, high thermal conductivity, fairly low cost, low health hazard,and high autoignition temperature (AIT). It is also beneficial to havefreezing point above the minimum start-up temperature, though thaw-outmeasures can be implemented. The AIT is also of minor importance, asinert or reducing-gas pressurization over the HTL would normally beincorporated; but it is still of some concern, should liquid leaksdevelop. Water, organic fluids, molten alloys, or molten salts willnormally be selected, based mostly on the temperature range. Table 1presents some pertinent data, some of which are estimates, on some HTLsat 500 K. The column labeled “Risks” gives a single, overall indicationof the three hazards normally considered—health, flammability, andreactivity.

The (turbulent-flow) tube-side heat transfer coefficient may becalculated by,

h _(tL) =B ₁ G ^(0.8) k _(t) ^(0.6)(C _(P)/μ)^(0.4) d ^(−1.8)  [3]

where μ is the dynamic viscosity (cP, centipoise, which is identical to1 mPa-s, or 0.001 kg/m-s) and B₁ is a dimensioned factor that is nearlyconstant over a wide range of conditions but varies with surfacefeatures and other exchanger design details. (Note: a fluid with μ=1 cPand ρ=1000 kg/m³ has kinematic viscosity, μ/ρ, of 1 cSt, centistokes.)Mew simple manipulations and calculations are useful:

G ^(0.8) =B ₂(ρν)_(0.8) d ^(1.6)  [4]

h _(tL) =B(ρν)^(0.8) k _(t) ^(0.6)(C _(P)/μ)^(0.4) d ^(−0.2)  [5]

F _(H)=ρ^(0.8) k _(t) ^(0.6)(C _(P)/μ)^(0.4)  [6]

h_(tL)=Bν^(0.8)F_(H)d^(−0.2)  [7]

where the B's are dimensioned constants, and F_(H) is a convenient,composite fluid property. A typical magnitude of B at a Reynolds numberof 10,000 to 20,000 inside smooth tubes is ˜5.6, assuming the parametersare in the units shown above. For 40 wt engine oil at 500 K, for v=10m/s in tubes of 0.0077 m ID (Re □ 15,000), this gives h_(tL) □19000W/m²-K. For comparison, FP Incropera gives a representative value ofoverall h_(t) for air in cross-flow with water inside finned tubes as˜35 W/m²-K, and maximum overall h_(t) of 6000 W/m²-K for steamcondensers.

From eq. 7 and the above example, it might appear that one simply needsto increase the HTL flow velocity to make h_(tL) very large compared tomean shell-side heat-transfer coefficient h_(S) (as desired for economicoptimization), but of course that consumes power—which increases almostas the cube of ν. The pumping power also increases with increasing ρ, μ,and flow length. Considering this, a better HTL figure of merit(composite fluid property) for its selection than the above F_(H) is thefollowing F_(M),

TABLE 1 HTL Properties at 500 K pour point, n.b.p. AIT, ρ, C_(P), kJ/kgμ, k_(t), W/m- F_(D) Name K K K kg/m³ K cP K Risks F_(H) F_(M) kDtacetone 185 329 738 411 3.44 0.05 0.08 2 146 1430 2200 ethanol 200 352636 475 3.43 0.06 0.09 1 164 1460 2400 butanol 210 380 699 581 3.780.094 0.10 1 179 1260 2300 water 274 373 — 835 4.57 0.11 0.646 0 7435130 22000 toluene 190 384 808 640 2.51 0.12 0.077 2 127 357 1030 cumene130 425 697 661 4.57 0.15 0.108 2 188 1090 2220 ethylene glycol 260 470673 935 3.16 0.34 0.2 1 221 369 1740 1-butylnaphthalene 260 561 800 8242.1 0.35 0.093 2 106 269 450 Delo 100 30 wt 243 570 550 670 2.5 0.3 0.090 100 300 500 PAO, Delo 400 5W40 230 580 620 670 2.5 0.3 0.09 0 100 300500 Delo 6170 40 wt 255 620 640 680 2.5 0.35 0.09 0 96 260 440 POE,Mobil 254 212 640 672 700 2.5 0.4 0.1 1 99 250 440 dioctyl phthalate 250657 780 798 2.1 0.5 0.11 1 98 200 365 1-dodecyl-naphthalene 305 676 800795 2.5 0.41 0.092 1 103 250 445 tri-o-cresyl phosphate 260 693 680 9502.2 0.4 0.11 2 136 340 680 TBPP-100 phosphate 270 708 795 900 2.2 0.50.13 1 123 255 515 polyphenyl ether 5P4E 280 749 860 970 1.9 0.6 0.13 1114 200 400 60NaNO₃—40KNO₃ 480 870 870 1950 1.4 4.5 0.45 2 166 77 27055Bi—45Pb 400 1800 — 10000 0.15 2.7 4 2 1150 300 2220 38Pb—37Bi—25Sn 4001900 — 9000 0.18 2.5 8 1 1770 520 5200

F_(M)=k^(0.6) _(t) (ρC_(p))^(0.8)/μ  [8]

The combination of the need to achieve h_(tL)>>h_(S) at moderate ν andflow length, along with good W matching, imposes constraints on the tubediameters and the tube paralleling scheme. The HTL would usually havenearly constant velocity throughout most of the cores, so tube diameterswould be nearly constant throughout. However, in cores containingseveral parallel tubes, it may be beneficial for them to combine attheir core entrances and exits to simplify tubing inter-connectionsbetween cores. Clearly, the HTL velocities in the interconnections couldbe very different from the typical value in the cores.

As seen in Table 1, F_(M) is low for organics compared to that for wateror molten alloys, but it is usually higher than that of molten salts—aconcept that has previously been misunderstood. Other advantages oforganics may include no freezing problems, no metal erosion, lowercorrosion, lower density, lower toxicity, lower cost, lower viscosity,and simpler disposal problems. Pressurized water can be used well beyond500 K, but exchanger costs are increased because of the very highstresses. An organic of low vapor pressure is often better, though insome applications lower-boiling fluids, such as ethanol or even acetone,could meet the specific requirements and be preferred. Note that therelative merits of the HTLs are temperature dependent.

Silicone fluids (such as Dow Corning 550, AIT of 755 K, but not suitablefor long term usage above 550 K) and low-grade hydrocarbon (HC)mixtures, such as Exxon Caloria HT-43 (AIT of 627 K) have been used.Some more attractive organic fluids with n.b.p. and AIT both above 660K, pour point below 320 K, and acceptable chemical stability and safetyare: (A) polyphenyl ethers (PPEs, aerospace lubricants anddiffusion-pump oils, 5-ring type 5P4E has AIT-880 K, n.b.p.=749 K, 290 Kpour point, ΔG_(f) ˜2 kJ/g, non toxic, has been used in short-termvapor-phase lubrication up to 870 K), (B) polyol esters (POEs, mosttype-2 aviation turbine oils, AIT usually ˜670 K, but AIT and n.b.p. canbe over 740 K), (C) polyalphaolefins (PAOs, a major component in type5W50 synthetic engine oil, 16 cSt at 100° C., AIT often ˜650 K, but AITcan be ˜700 K in heavy PAOs), (D) phosphate esters (used in aviationhydraulic fluids), (E) phenyl silicones, (F) fluorocarbons, (G) polymeresters (PEs), (H) phthalates, and (I) mixtures of the above andhigh-boiling alkylated polynuclear aromatics. See Table 1 for data ontwo alkylated polynuclear aromatics.

Highly branched alkanes are preferred to n-alkanes in engine lubricationapplications, as they have much better oxidation resistance, much lowerviscosity for a given boiling point, and are more resistant todehydrogenation and cracking. The relative price of such synthetic oils,similar to PAOs, should drop substantially over the coming decade.

Inexpensive tin-lead alloys may be acceptable as an HTL at hightemperatures with stainless or superalloy tubing. The solubility of ironin tin is about 0.1% at 650 K, and this may lead to excessive exchangererosion with low-alloy steel tubing (even after the molten alloy becomessaturated with iron, as there will be some thermal gradients in theliquid). The solubility of iron in both bismuth and lead is at least anorder of magnitude lower than in tin. However, alloys of more than 50%bismuth expand upon freezing (if not immediately, then after severaldays), and this could produce unacceptably high stresses within theexchangers. Lead-bismuth-tin alloys of relatively low tin content shouldbe fine with some low-cost steel alloys for the tubing. The38Pb-37Bi-25Sn alloy shown in Table 1 has an excellent balance of lowiron solubility, low vapor pressure, low toxicity, high F_(M), low cost,and low liquidus temperature, though perhaps lower Bi and Sn contentswith increased Pb and minor additions of antimony (Sb) would give aneven better balance.

Molten salts, especially mixtures of NaNO₃, KNOB, NaNO₂, and Ca(NO₃)₂,have often been used for HTLs. Some have freezing points lower thanthose of some lead alloys, but their upper temperature limits are lower.For example, KNO₃decomposes at 670 K and NaNO₃ decomposes at 650 K,though some mixtures, such as the eutectic listed in Table 1, havehigher stable temperature limits. There are some security risks, as allcan easily be used to make powerful explosives of limited stability.Moreover, their NEPA health ratings are usually “2, highly hazardous”,and their F_(M) is quite inferior to other options. When hot, they reactvigorously with most pump lubricants and elastomeric seals, and theyslowly attack many alloys at grain boundaries. Another complication withsalts, alloys, and heavy polynuclear aromatics is that they are solid atroom temperature.

The thermal conductivities of the gases expected in some emergingapplications typically range from 0.04 to 0.06 W/m-K at 500 K (for CO,C₄H₁₀, air, and some H₂/CO₂ mixtures of interest), gas densities areoften ˜5 kg/m³, C_(P) is often 1 to 3 kJ/kg-K, and μ is typically0.01-0.03 cP. For gases under the relevant shell-side (substantiallylaminar) conditions, a composite fluid property more useful than F_(M)for estimating the ease with which heat transfer may be achieved is

F _(G) =k _(t)(ρC _(P))²/μ  [9]

A useful expression for comparing liquid and gas heat transfer fluidsfor similar flow geometries (same hydraulic diameters, flow lengths,etc.) is (F_(M)F_(G))^(0.5), and this suggests the heat-transferchallenges for the gases could be two to three orders of magnitudegreater than for liquids for similar geometries. A simpler parameter,F_(D), for comparing diverse fluids is included in Table 1 and discussedin the last section with reference to applications for shell-sideliquids. The important point here is that there is usually little needto worry about tube-side heat-transfer enhancement. This allows enormousmanufacturing simplifications. The focus needs to be primarily onreducing passage thickness and increasing surface area on theshell-side.

FL Module Implementations. As noted previously, the core sets will beinside a pressure vessel sufficient for the shell-side pressure. Often,all the sets associated with the first shell-side stream (usually a gas)would be inside one pressure vessel, and those associated with thesecond shell-side stream would be inside a second pressure vessel. Thepressure vessel with the cores it contains is referred to as afluid-to-liquid (FL) recuperator module, and a typical embodiment isshown crudely in FIG. 3.

A typical FL recuperator module might contain 30 series-connected,thermally isolated finned-tube cores 31 (though the figure shows just 8cores for better clarity), each having typical external dimensions ofabout 1 m×1 m×0.03 m. The shell-side entrance and exit ports 32, 33 arenormally at the opposite ends from the tube-side entrance and exitports, 34, 35. A typical core is better illustrated in FIG. 4, thoughagain not likely to scale. Each 1 m×1 m core might typically have 40parallel finned tubes 41, each of 8 mm ID and 10 mm OD, each traversingthe full width, with typical center-to-center spacing of 25 mm, andtube-side entrance and exit manifolds 42, 43. FIG. 4, on the other hand,shows 20 tubes and 64 fins, which is closer to being typical for a 30cm×30 cm core, though even there the number of fins would likely begreater than shown by a factor of 2 to 4.

If the fin thickness is 0.5 mm and the fin pitch 1.6 mm, then theexample shell-side flow-section area A_(S) is about 0.7 m² and thetube-side flow-section area A_(T) is ˜0.002 m². Hence, A_(S) is ˜350times A_(T). For the typical core dimensions noted earlier (1 m×1 m×0.03m), the mean shell-side flow length L_(S) is about 0.03 m per core, andthe mean tube-side flow-length L_(T) is 1 m per core. Hence, L_(T) isabout 30 times L_(S). Note that this ratio is independent of the numberof cores when they are serially connected, as both flow lengths increaseby the same factor. The flow-section area ratio is also independent ofthe number of serially connected cores. There may be a substantial gapbetween adjacent cores, as depicted in FIG. 3, in the shell-side flowdirection for pressure equalization across the face of the cores andsome transverse mixing, but the shell-side flow is substantially axialwith respect to the pressure vessel.

The tube-side HTL flow is shown in FIG. 3 as being ducted 36 from anexit manifold on one side of a core diagonally across to the entrancemanifold on the next core. Note that the HTL enters all the cores on thesame side and exits all the cores on the opposite side. The diagonal HTLducting pattern is one way to improve tube-side flow homogeneity. Othermeasures may also be taken, and often a primary measure will bejudicious selection of the diameter of the tubes 41 such that the flowvelocity within them will achieve pressure drops that are large comparedto the pressure drop in the manifold while simultaneously meeting theother previously noted requirements with respect to pumping power,h_(t), and W_(L). Support structure for the cores is not shown, thoughclearly some is needed. The flow-cage 37 for constraining the shell-sideflow within the cores is only partially shown. For the conditionsnormally addressed, the shell-side volumetric flow rate will generallybe relatively high (especially when compared to tube-side), soshell-side pressure drops from inlet 32 to outlet 33 must necessarily below (to achieve low pumping power) and differential stresses on the cagecan easily be handled. The pressure vessel would preferably have a burstpressure greater than twice the mean shell-side relative pressure andgenerally much greater than 0.3 MPa.

One may define an HTL conductance Y_(F) [W] between adjacent cores as

Y_(F)=T_(d)W_(L)  [10]

where T_(d) is the mean temperature difference between adjacent cores.Adjacent cores are herein considered to be effectively thermallyisolated if the heat conducted between cores through solid materials isless than one-third of Y_(F). Such a condition is not easily met if morethan 20% of the fins are continuous from one core to the adjacent corein the shell-side flow direction, but such a condition is easily met ifnone of the fins are continuous between adjacent cores and the tubepattern is not interleaved between adjacent cores. However, adequatethermal isolation will sometimes be possible if up to 30% of the finsare largely continuous, except for holes for transverse pressureequalization, between adjacent rows.

It is not necessary for all tube rows to be thermally isolated. A“compound core” may have several rows of finned tubes thermally coupledby continuous fins between them for improved core robustness. However,the practical effectiveness limit is strongly dependent on the totalnumber of thermally isolated cores in series. Hence, it will often bedesirable for this number to be more than 20, though there will be somecases when as few as two thermally isolated cores per FL module aresufficient. It is unlikely that a compound core would contain more thanfour rows of thermally coupled rows of finned tubes. In most cases, eachthermally isolated core would be a single row of finned tubes, as shownin FIGS. 3 and 4.

Shell-side flow homogeneity is also essential for high ε, at least whenthe W's are similar. In most cases, allowing for pressure equalizationacross the faces, as is readily achieved when none of the fins arecontinuous between adjacent cores, will be sufficient, as shown in FIG.3. In the prior art, all the fins are usually continuous betweenadjacent rows, such as seen in the Armstrong Duralite™ Plate Fin Coilsproducts. A minor fraction could still be continuous. In such a case,pressure equalization across the faces of thermally isolated cores canreadily be achieved if holes or cut-outs are included in fins joiningadjacent cores.

With series-connected cross-flow exchangers, the flow homogeneity may befurther improved by inserting turbulent mixers in the gas flow streamsbetween cores. (This obviates the benefit from the flow routingillustrated in FIG. 2, but is better than the alternative ofchanneling—where, because of the viscosity dependence on temperature,the shell-side velocity may become higher than mean on one side of allthe cores when the shell-side gas is being heated a large amount in eachcore.) The use of separate, series-connected FL modules furthersimplifies the insertion of turbulent mixers into the shell-side stream.

FIG. 5 illustrates a portion of a core with a serpentine pattern thatmay be desired to better meet the HTL velocity and pressure-dropobjectives in some cases. If the tube-side flow for the core of FIG. 4were instead handled by 10 parallel tubes of 8 mm ID, each traversingthe full width 5 times in a serpentine pattern with center-to-centerspacing of 20 mm, the tube-side flow-section area then would be 5E-4 m².With a reasonable allowance for the bends at each end, each tube maythen need to be ˜6 m long. In this case, L_(T) would be about 200 timesL_(S), and A_(S) would be ˜1400 times A_(T). As in FIG. 4, theshell-side flow in FIG. 5 is normal to the plane of these figures.

A method of arranging thermally isolated, serially connected finnedtubes without manifolds between them is shown in FIG. 6. Such anarrangement may have manufacturing advantages in some cases. An optionis to stack a large number of serpentine finned tubes normal to theplane of FIG. 6, with the shell-side flow as indicated. For theshell-side flow direction assumed in FIGS. 4 and 5, the fins wouldnormally be continuous between the tubes as shown within a core.However, the fins could not be continuous between the tubes for the flowdirection shown in FIG. 6 and achieve thermal isolation between theseserially connected tubes. Conventional usage may not refer to such anarrangement of a single, serpentine finned tube as a “core”. When alarge number of the serpentine finned tubes as shown in FIG. 6 arestacked normal to the plane of the Figure, five thermally isolatedplanar cores are effectively formed.

As discussed earlier, it will sometimes be preferable to utilize morethan one liquid loop. Hence, in some applications, there may be two oreven three liquid loops servicing the cores in a single FL module. Insome applications, it may be preferable to utilize separate pressurevessels for the high-temperature cores, mid-temperature cores, andlow-temperature cores, and in such cases in particular a small number ofthermally isolated cores per FL module may be sufficient. In largerapplications, it will often be desirable to arrange modules in parallel,as it may not be optimum to manufacture modules larger than can easilybe transported by truck. Note that paralleling also does not affect theratios A_(S)/A_(T) or L_(T)/L_(S), but the ratio A_(S)/L_(S) steadilyincreases with capacity in an optimum design.

For very large modules, a hexagonal arrangement of the cores as shown inFIG. 7 (depicted without the containment vessel) may be preferred, as itpermits a larger ratio of A_(S)/L_(S) within practical truckingconstraints. Here, the shell-side fluid flow is generally radial, andthe core arrangement of FIG. 6 is assumed, though the arrangement ofFIG. 4 could also be used. Similar arrangements of finned-tube cores,except square rather than hexagonal, are commonplace in the AC industry,where shell-side air flow through a condenser exhausts to atmosphere.However, the prior-art condensers (A) utilize tube-side phase change formost of the enthalpy change, (B) are not enclosed in a pressure vessel,and (C) may not include serially connected thermally isolated cores.

For the hexagonal arrangement shown in FIG. 7, the shell-side face flowsare all functionally in parallel. Hence, the tube-side flows must alsobe functionally in parallel. In other words, all of the innermost coreswould connect to the same HTL port, and all of the outmost cores wouldconnect to the same HTL port. Obviously, pentagonal, octagonal, or othercircumferential arrangements of cores could also work well. The pressurevessel would normally have its axis aligned perpendicular to theshell-side flows through the cores.

Cores of significantly different characteristics may also be combined,either in series or in parallel, with predictable results, though theanalysis is more complex. Clearly, many variations in dimensions andpatterns are possible, but generally A_(S) would be more than 100 timesA_(T) and L_(T) would be more than 10 times L_(S). Such ratios appear tobe well outside the prior art in multi-pass, finned-tube, shell-and-tubeheat exchangers.

Core Modifications for Severe Conditions. For high performance atdemanding conditions (high temperatures, oxidizing atmospheres, or largedifferences in pressures between the two gases), appropriate changes inchoice of materials for the tubing, fins, and braze are required. Thetubing material is selected primarily for yield strength at the requiredtemperature, formability, brazability, and corrosion resistance. The finmaterial is selected primarily for thermal conductivity, cost, corrosionresistance, melting point, and brazability. In some cases, the fins havebeen pressed on rather than brazed on, but this approach is lessdesirable for extremes of temperature, for closely spaced thin fins, orif much vibration is likely.

Alumina-dispersion-strengthened copper, aluminum, or nickel areparticularly good choices for the fins, though cobalt and alloys arealso possible for high temperature fins. While most superalloys havepoor thermal conductivity compared to pure metals near room temperature,some with superior oxidation and corrosion resistance, such as Haynes214 (16Cr, 4.5AL, 3Fe, 0.2Y, bal-Ni), have fairly good thermalconductivity at high temperatures (32.4 W/m-K at 1255 K).

Some superalloys, such as Haynes 188 (38Co, 22Cr, 22Ni, 14.5W, 2Fe, 1Mn,0.3Si, 0.1C, 0.07La), have good brazability and formability in theannealed state as well as outstanding oxidation resistance andhigh-temperature strength (˜1400 K for 70 MPa yield strength in alloy188). An alloy similar to Haynes 188 would be well suited forhigh-temperature exchanger tubing, though modifications to reduce costand improve formability and brazability, particularly by reducing Co, W,and Cr, may be preferred. The tubing material can have poor thermalconductivity with little consequence on performance. If the hydrostaticpressure on the HTL(s) is maintained near the mean of the pressures ofthe gas streams, preferably within a factor of two of this mean, thestresses on the tubing are reduced.

Brazes compatible with the higher temperatures and materials arerequired. Nickel-plated dispersion-strengthened-copper fins could bebrazed to such using filler BNi-7 (890° C. liquidus, 85Ni, 14Cr, 10P).Superalloy or alumina-dispersion-strengthened nickel fins could bebrazed to Haynes 188 or similar tubing using BNi-5 (1135° C. liquidus,70Ni, 19Cr, 10Si) for operation at still higher temperatures. Methodsfor applying chromium platings to the finned tube rows can be developed,based on the prior art.

Organic HTLs at High Temperatures. An organic HTL may be used quitesatisfactorily at a much higher temperature than that for which it hasnormally been recommended if suitable measures are taken. First of all,it is most important that the surfaces in contact with the hot oil(tubing interiors, etc.,) be catalytically deactivated with a thin layer(0.1 micron is sufficient) of coke—carbon and very heavy condensedpolynuclear aromatics. Thermal (non-catalytic) reactions require muchhigher temperatures than catalytic, and most metallic or oxide surfaceshave some catalytic activity. Secondly, since water catalyzes reactionson many metal surfaces, it is important to maintain the liquid pressurewell above the maximum external gas pressure (the greater of thepressures in shell-side gas-1, gas-2, and ambient) at all times toprevent ingress of air and moisture through minute leaks. Of course, itis important to insure that any organic HTL is initially de-gassed ofdissolved O₂ and H₂O.

In general, there are four primary types of thermal reactions that willdominate for most of the heavy HCs likely to be used for ahigh-temperature HTL: cracking, dehydrogenation, de-isomerization, andaromatic polymerization or condensation. All but de-isomerization(conversion from a highly branched to a less branched structure) aresomewhat inhibited by moderate H₂ and CH₄ concentrations within theHTL—or perhaps it is more proper to say that high H₂ and CH₄concentrations increase the rate of the reverse of many undesiredreactions.

As previously noted, a small surge tank is required for the HTL toaccommodate expansion and contraction. To extend the lifetime anduseable temperature limit, the gas overhead 7 in this reservoir shouldhave an H₂ partial pressures of at least 0.01 MPa and possibly as muchas 5 MPa, though excessive H₂ partial pressures will increase cracking(especially of n-alkanes) and hydrogenation of aromatics intolower-boiling cyclics. Hence, it may be desirable to also havesignificant methane partial pressure, possibly as much as 15 MPa, as itis less reactive. For some HTLs, such as water, glycols, phthalates,silicones, polyol esters, and polyphenyl ethers, partial pressurizationwith argon and perhaps N2 may be preferred. Maintaining an excessivetotal pressure on the HTL increases the cost of the high-temperaturecores and exacerbates problems with dynamic seals, but an HTL staticpressure about 0.1 to 1 MPa above the higher of the shell-side gaspressures would normally add little to the system cost.

The concentrations of CH₄ and H₂ dissolved within the HTL are determinedlargely by their partial pressures and the liquid temperature in the HTLreservoir. Solubilities of H₂ in HCs (A) are generally higher foralkanes than for aromatics, (B) they increase with increasingtemperature, (C) they approximate a Henry's law behavior, and (D) theydecrease slowly with increasing molecular mass of the HTL. Thesolubility of H₂, in moles H₂ per kg liquid per MPa, at 460 K are 0.068and 0.044 for hexadecane (C₁₆H₃₄) and tetralin (C₁₀H₁₂) respectively,for example. Solubilities at 520 K are about 30% higher. Solubilities invery heavy oils are about half that for hexadecane. Methane solubilityis much higher (by perhaps a factor of 20 at 460 K) and much lessdependent on temperature. When the HTL cools during power-down, it mayeffervesce H₂.

It will normally be preferable to have the liquid pumps at thelow-temperature points in the loops, as shown in FIG. 1, as thissimplifies problems associated with dynamic seals. It may also bepreferable to have the reservoir near the low-temperature point in theloop to avoid H₂ super-saturation within the HTL at its cooler points inthe loop, as super-saturation could lead to hydrogen effervescence inthe cooler exchangers and reduced heat transfer. However, some level ofH₂ super-saturation is usually quite stable in HCs, and this may furtherinhibit production of coke precursors with some HTLs. Hence, it may bepreferable to have the surge tank at a higher temperature point in theloop, even though this increases its cost a little.

Even with the above measures, operating at temperatures near the upperpractical limits will result in the production of reaction products,both light and heavy, that are undesirable beyond a certain level butare quite tolerable at low levels. In most cases, this will simply meanthat periodic HTL changes will be required. For large installationsthere are other options. Cracking produces light alkenes, some of whichwill be hydrogenated to light gases such as C₂H₆, C₃H₈, and C₄H₁₀, whichare less than optimum for reservoir pressurization. An easy way to dealwith such is to continually, slowly vent some pressurization gas andmaintain the desired pressure with fresh gas of optimum mixture. Ofcourse, membranes and other separations methods could be used toseparate the vented gas into useful product streams if desired. Some ofthe alkenes will alkylate with other alkanes or aromatics to heavy HCsand coke precursors in the HTL. One way to maintain the HTL at anacceptable composition is to steadily bleed HTL from the reservoir andmaintain the desired level with fresh supply. Various separationsmethods could be applied to the used fluid for reclamation. Moreexamples of reaction-product separation processes are disclosed in aco-pending patent application on Dual-source Organic Rankine Cycles.

In summary, the following are required to operate with organics at hightemperatures:

1. Deactivate all surfaces in contact with the hot HTL.2. Maintain sufficient HTL pressure to prevent ingress of air andmoisture through minute leaks.3. Maintain an optimum gas mixture pressurizing the HTL.4. Remove primary HTL reaction products before they lead to excessivecoking.5. Select a fluid with high chemical stability with optimum gaspressurization.

For the temperatures indicated in the example of FIG. 1 with appropriategas pressurization over the HTLs, the HTL for sets A and B could bedioctyl phthalate, a PAO oil, or a POE oil. For sets C and D, a moltenalloy, a molten salt, PPE-5P4E, or possibly an alkylated polynucleararomatic could be used.

Cryogenic Applications. While some of the largest applications may be atelevated temperatures in chemical processes and power plants, there willalso likely be enormous applications at cryogenic temperatures, as veryhigh ε in exchange between gases is often required there. Moreover, gasviscosities (and hence pressure drops) there are often so low that it isvery difficult to establish the uniform flow conditions that areessential for high ε. As previously noted, the use of separate,series-connected cores or FL modules allows simple insertion ofturbulent mixers into the shell-side stream between modules.

For cryogenic applications, the fin pitch can be further reduced—because(A) viscous losses are much smaller, (B) the fin-metal thermalconductivity can be an order of magnitude higher, (C) the gas thermalconductivity is often much lower, (D) the HTL may have a higher F_(M),and (E) corrosion is more readily controlled.

One HTL, cumene (isopropylbenzene, C₉H₁₂), is listed in Table 1 that isparticularly advantageous down to 130 K, and others are suitable forlower temperatures. Propane, for example is usable down to 90 K and iseasily liquefied at room temperature, as its critical temperature T_(C)is 370 K. For lower temperatures, gases with T_(C) well below 300 K arerequired, and this complicates start-up somewhat, as a rather largecompressed-gas reservoir is needed. Oxygen (T_(C)=155 K) is an excellentHTL for the range of 60-130 K. Fluorine oxide, F₂O (T_(C)=215 K), issuitable for the 55-170 K range, and other gases can be used over other,narrow ranges. For example, H₂ (T_(c)=33 K) can be used over the 15-30 Krange. However, very high pressures are required to condense these gasesnear the upper ends of their maximum liquid ranges, and this increasesexchanger cost.

In principle, a gas could be used as the heat transfer intermediary,where the obvious choice for the 35-60 K range would be hydrogen.However, high (tube-side) h_(tL) with low pumping power cannot beachieved with a gas as the intermediary above its T_(C), as its densityis much too low at practical pressures. The best way to increase h_(tL)with gases is to use MMP tubing, which indeed works beautifully at veryhigh pressures.

The minimum competitive size of the compound exchanger for cryogenicapplications will be smaller than for most high-temperatureapplications—because mass is often much more critical and meantemperature difference between the gas streams may need to be an orderof magnitude smaller. The compound exchanger seems likely to bepreferred in many cryogenic recuperators down to 90 K at exchange powersabove 1 kW for gas pressures below 0.5 MPa.

Compact Recuperator Variations. An advantage not yet noted for thecompound exchanger is that it can greatly reduce the ducting costs inlarge plants where the heat generated in one process is needed inanother process hundreds of meters or even tens of kilometers away. Insuch a case, it will sometimes be easier to achieve optimum thermalbalancing—matching the gas W and temperature to those of the HTL withineach module—by splitting and re-combining the HTL streams at numerouspoints. When streams are combined, their temperatures should be similarfor minimal exergy destruction. A portion of an HTL may be split outfrom an intermediate point in an exchanger module to exchange energywith another process and then be re-combined at an appropriate pointwhere the temperature is similar.

Of course, it is not uncommon to transfer heat long distances usingeither phase change (usually water) or liquids (including many of thosementioned earlier as good HTLs for compound recuperators). For example,Severinsky in US publication 2006/0211777 notes that it can beadvantageous to transfer heat throughout a large plant using a number ofdifferent phase-change heat-transfer fluids (HTFs).

While it is important to emphasize that exergy destruction is morereadily minimized by avoiding substantial phase change when there is alarge temperature difference between the hot source gas and the coldsource gas (so the number of HTL loops can be reduced), a minor amountof boiling and condensing may take place within the HTLs withoutdeparting from the spirit of this invention. Hence, the HTL may bereferred to as an HTF, as is customary in the prior art, though in ahigh-ε recuperator the enthalpy associated with phase change would besmall compared to that associated with temperature change.

The description of the shell-side fluids as “clean gases” in earlierdiscussions requires further clarification. It is anticipated that inmany cases the amount of condensation, acid formation, ice formation,corrosion, and particulates will be minor, though such are notprecluded. When fouling mechanisms are negligible, the fin pitch may bereduced for improved compactness. However, the FL recuperator will stillbe advantageous in many applications where these mechanisms aresubstantial—though perhaps not where they are strongly dominant.

Fouling will often be significant in only one of the gas streams, andoften only at either the hot or cold end of that stream. A strongadvantage of the compound recuperator is that it may readily permitindividual modules to be switched off-line for rejuvenation (defrosting,cleaning, re-plating, etc.) while a fresh module is put into service. Insome case, the fouled module may need to be shipped back to the factoryfor service, but often it will simply need to be drained, defrosted,burned out, or solvent washed. In many cases, it will simply benecessary to orient those modules in which significant condensationoccurs so that the condensate readily drains while in use—as for examplethe draining of moisture from the common AC evaporator on a humid day.The compound exchanger will often permit a dramatic reduction in thenumber of replacement exchanger modules that will need to be kept onhand in a large process plant.

Large applications are also anticipated where the shell-side fluids areviscous organic liquids, since such exchanges also benefit from the veryshort flow passages that can more easily be obtained in the inventivemodule. While the benefits may be greatest with oils of high viscosity,even moderate-viscosity oils, such as 1,3-diphenylpropane at 310 K,where μ=4.4 cP, k_(t)=0.12 W/m-K, ρ=968 kg/m³, and C_(P)=2 kJ/kg-K,would benefit when high effectiveness is needed, especially ifeffervescence is also present in one of the streams. In such a case, aphase separator or flash drum can be inserted between modules or evenbetween cores to separate the evolved shell-side gas so the fluid'svolumetric flow rate (and hence velocity) remains low—to limit viscouslosses.

A composite fluid property that is dimensionally much simpler than(F_(M)F_(G))^(0.5) and nearly as valid for comparing diverse fluids forsimilar flow geometries is

F _(D) =k _(t) ρC _(P)/μ.  [11]

The HTL in Table 1 having the lowest F_(D) (i.e., least desirability) at500 K is (again) the salt, where F_(D)=2.7E5 J²/(s-m⁴-K²-cP) in thesemixed reduced units, which herein will be abbreviated Dt (for Doty). (InSI units, 1 Dt=1000 J²/(kg-m³-K²).) For comparison (again at 500 K),F_(D) is seen to be ˜440 kDt for 40 wt engine oil and 22 MDt for water.

In contrast, the shell-side fluids have lower F_(D). A typical value forthe gas conditions indicated earlier (500 K, 5 kg/m³, 0.05 W/m-K, etc.)would be ˜25 kDt. Some liquids for which high-performance heat recoverywill be needed have F_(D) well below those of a preferred HTL, and insuch cases, a compound recuperator can be advantageous, particularly ifthe temperature permits the use of a tube-side HTL of very high F_(D),such as water or a molten alloy.

For diphenylmethane at 310 K for example, F_(D) is ˜100 kDt, and for1,3-diphenylpropane F_(D) is 52 kDt. For heavy oils, F_(D) can be anorder of magnitude smaller yet, even at temperatures where substantialheat recuperation may be needed in some situations.

The FL recuperator will be useful for heat recovery in many fluids whereF_(D) is less than 200 kDt at the operating conditions, which generallyimplies μ≧1 cP for organic liquids. When the shell-side fluid has ratherhigh F_(D) (as for some low-viscosity liquids and gases at very highpressures), a tube-side HTF would be needed with very high F_(D), suchas water or a molten alloy. However, a tube-side HTF with F_(D) as lowas 200 kDt would be satisfactory when operating with shell-side fluidsof rather low F_(D). Preferably, the tube-side fluid would have F_(D)more than 10 times that of the shell-side fluids (which of course can bevery different, and at very different conditions).

Although this invention has been described herein with reference tospecific embodiments, it will be recognized that changes andmodifications may be made without departing from the spirit of thepresent invention. All such modifications and changes are intended to beincluded within the scope of the following claims.

1. A fluid-liquid (FL) recuperator for heat exchange between ashell-side fluid stream of mean flow-section area A_(S) and a tube-sideheat-transfer liquid (HTL) of mean flow-section area A_(T), saidrecuperator comprising a plurality of thermally isolated seriallyconnected adjacent exchanger cores, wherein cores are consideredthermally isolated if fewer than 30% of the fins are substantiallycontinuous in the shell-side flow direction between adjacent cores andthe tube pattern is not interleaved between adjacent cores, said corefurther characterized as comprising at least one substantially planarrow of finned tubes for transfer of heat between a shell-side stream ofmean flow length L_(S) and an HTL of mean tube-side flow length L_(T),said core further characterized in that the length of the fins per rowin the shell-side flow direction is less than 80 mm, the fin pitch isless than 8 mm, A_(S) is greater than 100A_(T), and L_(T) is greaterthan 10L_(S), a cylindrical pressure vessel having burst pressuregreater than 0.3 MPa enclosing said cores, aligned such that its axis issubstantially parallel to the shell-side flow, and having shell-sidefluid-flow ports near its opposite ends, means for directing theshell-side flow substantially through said cores, said FL recuperatorfurther characterized in that the fluid flow connections between saidcores are such as to achieve a substantially counterflow exchangebetween the shell-side stream and the HTL.
 2. The FL recuperator of 1 inwhich said cores are further characterized in that the length of thefins per row in the shell-side flow direction is less than 40 mm, thefin pitch is less than 4 mm, A_(S) is greater than 200A_(T), and L_(T)is greater than 20L_(S).
 3. The FL recuperator of 1 furthercharacterized as including a transverse passage between thermallyisolated cores to equilibrate shell-side pressures across the faces ofsaid cores.
 4. The recuperator of 1 further characterized as utilizingalumina-dispersion-strengthened-metal fins brazed perpendicularly ontoalloy tubes of yield strength greater than 70 MPa at 750 K.
 5. Therecuperator of 1 further characterized as utilizingalumina-dispersion-strengthened-nickel fins brazed perpendicularly ontosuperalloy tubes of yield strength greater than 70 MPa at 1300 K.
 6. TheFL recuperator of 1 in which said tubes are further characterized ashaving a substantially smooth inner surface with a coating for catalyticdeactivation.
 7. The FL recuperator of 1 further characterized ascomprising a minimum of five thermally isolated cores.
 8. The FLrecuperator of 1 further characterized as including multiple pairs ofports for multiple liquid loops.
 9. A fluid-liquid (FL) recuperator forheat exchange between a shell-side fluid stream of mean flow-sectionarea A_(S) and a tube-side heat-transfer liquid (HTL) of meanflow-section area A_(T), said recuperator comprising a circumferentialarrangement of a plurality of thermally isolated serially connectedadjacent exchanger cores, wherein cores are considered thermallyisolated if fewer than 30% of the fins are continuous between adjacentcores in the shell-side flow direction and the tube pattern is notinterleaved between adjacent cores, said core further characterized ascomprising a substantially planar row of finned tubes for transfer ofheat between a shell-side stream of mean flow length L_(S) and an HTL ofmean tube-side flow length L_(T), said core further characterized inthat the length of the fins per row in the shell-side flow direction isless than 80 mm, the fin pitch is less than 8 mm, A_(S) is greater than100A_(T), and L_(T) is greater than 10L_(S), a cylindrical pressurevessel having burst pressure greater than 0.3 MPa enclosing said coresand aligned such that its axis is substantially perpendicular to theshell-side flow through said cores, said shell-side flow beingsubstantially radial with respect to said pressure vessel, means fordirecting the shell-side flow substantially through said cores, said FLrecuperator further characterized in that the fluid flow connectionsbetween said cores are such as to achieve a substantially counterflowexchange between the shell-side stream and the HTL.
 10. The FLrecuperator of 9 further characterized as including transverse passagesbetween thermally isolated cores to equilibrate shell-side pressuresacross the faces of said cores.
 11. A method for heat exchange between afirst shell-side fluid stream at mean pressure p₁ and a secondshell-side fluid stream at mean pressure p₂, said method using a firstset of serially connected thermally isolated cross-flow exchanger coresfor transfer of heat between an intermediary tube-side heat transferfluid (HTF) and the first shell-side stream, a second set of seriallyconnected thermally isolated cross-flow exchanger cores for transfer ofheat between the HTF and the second shell-side stream, said HTFcharacterized as being substantially liquid phase throughout all coresand having critical temperature not less than 370 K, wherein a core ischaracterized as comprising at least one row of finned tubes, saidfinned tubes are further characterized in that the length of the tubefins per row in the shell-side flow direction is typically less than 80mm and the fin pitch is typically less than 8 mm.
 12. The method of 11further characterized as having more than 4 thermally isolated coresexchanging with each shell-side stream and having effectiveness εgreater than 60% at design operating conditions.
 13. The method of 11 inwhich said shell-side fluids are further characterized as selected fromthe set comprised of organic liquids having viscosity greater than 1 cPat 310 K and gases at pressure greater than 0.05 MPa.
 14. The method of11 where said HTF is further characterized as having flow rate G_(L)kg/s, specific heat C_(PL) J/kg-K, and W_(L)=G_(L)C_(PL), said firstshell-side fluid has flow rate G₁, specific heat C_(P1), andW₁=G₁C_(P1), said second shell-side fluid has flow rate G₂, specificheat C_(P2), and W₂=G₂C_(P2), said geometric mean shell-side conditionsdefined by W_(S)=(W₁W₂)^(0.5), said tube-side conditions furthercharacterized in that W_(L)>0.7W_(S) and W_(L)<1.4W_(S).
 15. The methodof 11 further characterized in that said HTF is selected from the setcomprised of water, organics, molten alloys; and molten salts and isfurther characterized as having F_(D) greater than 2E5 J²/(s-m⁴-K²-cP)at the mean operating temperature, whereF _(D) =k _(t) ρC _(P)/μ, where k_(t) is in W/m-K, ρC_(P) is in J/m³-K,and μ is in cP.
 16. The method of 15 further characterized in that saidtube-side HTF has F_(D) greater than the lesser of the F_(D) of eitherof the shell-side streams by more than a factor of 10 at mean operatingconditions.
 17. The method of 15 further characterized in that each ofsaid shell-side streams has F_(D) less than 2E5 J²(s-m⁴-K²-cP) at theoperating conditions.
 18. The method of 11 further characterized asincluding a plurality of liquid pumps and liquid reservoirs forcirculation of a plurality of HTFs.
 19. The method of 11 in which saidHTF is further characterized as substantially selected from the setcomprised of polyphenyl ethers, polyol esters, polyalphaolefins,phosphate esters, phthalates, silicones, fluorocarbons, polymer esters,organic liquid mixtures that include alkylated polynuclear aromatics,and engine oils.
 20. The method of 19 further characterized as includinga liquid reservoir with overhead gas space, said gas having H₂ partialpressure greater than 0.01 MPa, O₂ partial pressure less than 1 kPa, H₂Opartial pressure less than 10 kPa, and total pressure greater than 0.15MPa.
 21. The method of 11 in which said HTF is further characterized assubstantially comprised of a lead-bismuth-tin alloy.
 22. The method of11 further characterized in that the mean pressure in said HTF isbetween 50% and 200% of the mean of p₁ and p₂.
 23. The method of 11wherein one of said shell-side fluids is further characterized as anorganic solvent containing a dissolved gas that effervesces when thefluid is heated, and means are included between cores for separating theeffervesced gas from the liquid.
 24. The method of 11 wherein one ofsaid shell-side fluids is further characterized as a gas containing avapor that condenses when the fluid is cooled, with means for drainingthe condensed liquid from a core.
 25. The method of 11 in which said HTFis further characterized as an organic liquid, and means are includedfor separation of reaction products from said HTF.
 26. The method of 11further characterized in that p₂ is greater than 3p₁ and the typical finpitch in said second set of cores is less than 70% of the typical finpitch in said first set of cores.
 27. The method of 11 furthercharacterized as including transverse passages between thermallyisolated cores to equilibrate shell-side pressures across the faces ofsaid cores, wherein cores are considered thermally isolated if fewerthan 30% of the fins are continuous between adjacent cores in theshell-side flow direction and the tube pattern is not interleavedbetween adjacent cores.